A high torque density motor for electric aircraft

Abstract

The application discloses a high-torque-density motor for an electric aircraft, adopts an inner rotor, a 54-slot 42-pole mode, a four-section rotor core skew slot magnetic steel radial insertion mode, has the performance of symmetrical magnetic motive force waveform, high winding coefficient, low torque ripple and low cogging torque, and no end ring current circulation; the motor adopts a centralized double winding, a shell cooling fin air cooling mode, has the characteristics of large torque, light weight, simple cooling mode, large safety margin, etc. The application can realize higher torque density and power density, and the heat is stable, controllable and not easy to burn.

Description

[Technical Field]

[0001] This invention belongs to the field of motor technology, and in particular relates to a high torque density electric motor for aircraft. [Background Technology]

[0002] With the rapid development of the low-altitude economy, electric aircraft are placing increasingly higher demands on the torque and power density of their propulsion motors. Compared to motors in new energy vehicles, which generally have a power density of no more than 3.5 kW / kg and a torque density of no more than 8 Nm / kg, the highest torque density required for electric aircraft motors has reached 50 Nm / kg, and the highest power density has reached 8.5 kW / kg. However, the liquid cooling system for motors is more complex, and if the liquid cooling system fails, the motor will burn out, leading to irreversible consequences.

[0003] Therefore, it is necessary to provide a new high torque density electric motor for aircraft to solve the above-mentioned technical problems. [Summary of the Invention]

[0004] The main objective of this invention is to provide a high torque density electric motor for aircraft, which achieves higher torque and power density while maintaining stable and controllable heat generation and being less prone to burnout.

[0005] The present invention achieves the above-mentioned objective through the following technical solution: a high torque density electric motor for an electric aircraft, comprising a stator and a rotor disposed inside the stator; the stator comprises a stator core and windings, the inner wall of the stator core having a plurality of stator slots arranged at equal angles along its circumference, each stator slot containing the windings; the rotor comprises a rotor core and a magnet assembly, the outer wall of the rotor core having a plurality of receiving cavities arranged at equal angles along its circumference, each receiving cavity containing a set of magnet assemblies;

[0006] The stator has 54 slots and the magnet assembly has 42, forming a 54-slot, 42-pole configuration.

[0007] The winding adopts a concentrated double winding form;

[0008] The stator core has an outer diameter D1 = 355 mm and an inner diameter D2 = 308.6 mm.

[0009] The outer diameter of the rotor core is D3 = 306 mm and the inner diameter is D4 = 286 mm.

[0010] The air gap L1 between the stator and the rotor is 1.3 mm;

[0011] The stator slot has a slot opening width L2 = 3.5 mm, a slot depth L3 = 18.7 mm, a slot tip angle θ1 = 28°, and the stator slot is a parallel slot with a parallel slot width L4 = 9.4 mm.

[0012] The rotor adopts a four-segment skewed pole treatment, with skewed pole treatment angles of -0.5626°, -0.1875°, 0.1875°, and 0.5625° respectively.

[0013] Furthermore, stator teeth are formed between two adjacent stator slots, and the air gap is formed between the inner wall of the stator teeth and the outer wall of the magnet assembly.

[0014] Furthermore, the main cross-sectional shape of the stator slot is rectangular, and an inwardly recessed slot is formed on the side near the inner wall of the stator core, with the inwardly recessed profile being a straight line; the slot tip angle is the inclination angle of the inwardly recessed profile.

[0015] Furthermore, the polar arc angle θ2 = 168°.

[0016] Furthermore, the magnets in the magnet assembly have a thickness L5 = 5.2 mm and are inserted into the receiving cavity in a radial insertion manner.

[0017] Furthermore, the magnet is an N52UH neodymium iron boron magnet with a remanence of 1.46T and a maximum operating temperature of 180℃.

[0018] Furthermore, the group of magnets forms a pole, and the pole includes 6 to 10 magnets.

[0019] Furthermore, the winding is made of 240℃ temperature-resistant, corona-resistant polyimide copper flat wire.

[0020] Furthermore, the stator slot fill factor is 89.4%.

[0021] Furthermore, the stator slot is provided with insulating paper that isolates the inner wall of the stator slot from the winding; the insulating paper is 0.25mm thick DuPont Nomex T410 aramid insulating paper; the gap of the winding in the stator slot is filled with insulating varnish, the insulating varnish being VX4201 unsaturated polyesterimide resin.

[0022] Furthermore, the stator core and the rotor core are made of Hperco50 iron-cobalt-vanadium alloy sheets with a thickness of 0.15mm, and the stack height of the stator core and the rotor core is 65mm, with a stacking factor of 0.97.

[0023] Furthermore, a heat dissipation component is provided on the outer periphery of the stator; the heat dissipation component includes a heat dissipation ring and a plurality of heat dissipation fins arranged in a ring at equal angles along the outer periphery of the heat dissipation ring; the heat dissipation ring is tightly wrapped around the outer periphery of the stator core.

[0024] Compared with existing technologies, the advantages of this invention's high torque density electric aircraft motor are as follows: Through optimized design in key aspects such as motor material selection, magnet arrangement, skewed pole treatment, and winding form, it is applied to electric propulsion aircraft. The use of an internal rotor, 54 slots, 42 poles, and a four-segment rotor core with radially inserted skewed magnets results in symmetrical magnetomotive force waveforms, high winding coefficients, low torque pulsation and cogging torque, and no end-circulating current. The motor employs a centralized dual-winding design and a shell heat sink air-cooling system, offering advantages such as high torque, light weight, simple cooling, and a large safety margin. The motor designed in this scheme achieves higher torque and power densities while maintaining stable and controllable heat generation and minimizing the risk of burnout. [Attached Image Description]

[0025] Figure 1 This is a schematic diagram of the cross-sectional structure of an embodiment of the present invention;

[0026] Figure 2 This is a schematic diagram of the structure of the winding in the stator slot in an embodiment of the present invention;

[0027] Figure 3 This is a partial structural diagram in an embodiment of the present invention;

[0028] Figure 4 This is a magnetomotive force waveform diagram of a motor using a 54-slot, 42-pole configuration according to an embodiment of the present invention.

[0029] Figure 5 This is a histogram of the winding coefficient distribution of a motor with a 54-slot, 42-pole configuration according to an embodiment of the present invention.

[0030] Figure 6 This is a BH curve diagram of the Hperco50 iron-cobalt-vanadium alloy material according to an embodiment of the present invention;

[0031] Figure 7 This is a diagram showing the remanence, coercivity, and squareness coefficient Hk / Hcj of the N52UH magnet in this embodiment of the invention.

[0032] Figure 8 This is an electromagnetic simulation data graph at peak power in an embodiment of the present invention;

[0033] Figure 9 This is a magnetic induction density cloud map at peak power in an embodiment of the present invention;

[0034] Figure 10 This is a diagram showing the magnetic induction intensity of each part under peak power in an embodiment of the present invention;

[0035] Figure 11 This is a loss distribution diagram obtained from simulation at peak power in an embodiment of the present invention;

[0036] Figure 12 This is a TS curve obtained from simulation at peak power in an embodiment of the present invention;

[0037] Figure 13 This is a MAP diagram of efficiency obtained from simulation at peak power in an embodiment of the present invention;

[0038] Figure 14 This is a thermal simulation result diagram at peak power in an embodiment of the present invention;

[0039] Figure 15 This is an electromagnetic simulation data diagram under rated power in an embodiment of the present invention;

[0040] Figure 16 This is a magnetic induction density cloud map under rated power in an embodiment of the present invention;

[0041] Figure 17 This is a diagram showing the magnetic induction intensity of each part under rated power in an embodiment of the present invention;

[0042] Figure 18 This is a loss distribution diagram obtained from simulation at rated power in an embodiment of the present invention;

[0043] Figure 19 This is a TS curve obtained from simulation at rated power in an embodiment of the present invention;

[0044] Figure 20 This is a MAP diagram of efficiency obtained from simulation at rated power in an embodiment of the present invention;

[0045] Figure 21 This is a thermal simulation result diagram under rated power in an embodiment of the present invention;

[0046] Figure 22 This is a simulation result data graph showing the maximum torque in an embodiment of the present invention;

[0047] Figure 23 This is a magnetic induction density cloud map under maximum torque in an embodiment of the present invention;

[0048] Figure 24 This is a magnetic induction density diagram of each part under the maximum torque in an embodiment of the present invention;

[0049] Figure 25-26 This is a simulation diagram of rotor strength in an embodiment of the present invention;

[0050] Figure 27 The above diagram shows the simulation results of cogging torque and torque pulsation in an embodiment of the present invention.

[0051] Figure 28 This is a 2D spectrum of the time-dominant stator radial harmonics according to an embodiment of the present invention;

[0052] Figure 29 This is a Campbell sound pressure field spectrum diagram according to an embodiment of the present invention;

[0053] Figure 30 - This is a diagram showing the stator radial stress data of an embodiment of the present invention;

[0054] Figures 31-33 The figures show the phase current curve, peak terminal voltage, and back electromotive force (EMF) curve of an embodiment of the present invention.

[0055] The numbers in the diagram represent:

[0056] 100 - High torque density electric motor for aircraft;

[0057] 1-Stator, 11-Stator core, 111-Stator slot, 112-Stator tooth, 113-Insulating paper, 114-Insulating varnish, 12-Winding; 2-Rotor, 21-Rotor core, 22-Magnet assembly; 3-Heat sink, 31-Heat ring, 32-Heat fin.

Detailed Implementation Methods

[0058] Example 1:

[0059] Please refer to Figures 1-3 This embodiment is a high torque density electric motor 100 for aircraft, which includes a stator 1, a rotor 2 disposed inside the stator 1, and a heat sink 3 disposed outside the stator 1.

[0060] The stator 1 includes a stator core 11 and windings 12. The inner wall of the stator core 11 has a plurality of stator slots 111 arranged at equal angles along its circumference, and each stator slot 111 contains a winding 12. Stator teeth 112 are formed between adjacent stator slots 111. The outer diameter D1 of the stator core 11 is 355 mm, and the inner diameter D2 of the stator core 11 is 308.6 mm.

[0061] The rotor 2 includes a rotor core 21 and a magnet assembly 22. The outer wall of the rotor core 21 has several accommodating cavities (not shown in the figure) arranged at equal angles along its circumference, and each accommodating cavity contains a set of magnet assemblies 22. The outer diameter D3 of the rotor core 21 is 306 mm, and the inner diameter D4 of the rotor core 21 is 286 mm.

[0062] The air gap L1 between stator 1 and rotor 2 is 1.3 mm, which is the distance between the outer wall of magnet assembly 22 and the inner wall of stator core 11. The main cross-sectional shape of stator slot 111 is rectangular, and an inwardly narrowed slot is formed on the side near the inner wall of stator core 11. The inwardly narrowed profile is a straight line, and the slot width L2 is 3.5 mm. The slot depth L3 is 18.7 mm, and the inclination angle of the inwardly narrowed profile, i.e., the slot tip angle θ1, is 28°. Stator slot 111 is a parallel slot, and its parallel slot width L4 is 9.4 mm. The pole arc angle θ2 is 168°, and the magnet thickness L5 is 5.2 mm. The magnet is 52UH neodymium iron boron magnet with a remanence of 1.46 T and a maximum operating temperature of 180℃.

[0063] The more stator slots 111 there are, the lower the motor's harmonic content, effectively reducing additional losses and harmonic leakage reactance, making the magnetomotive force waveform closer to a sine wave, which is beneficial for increasing motor torque. Additionally, the increased total heat dissipation area around the coils in the slots facilitates heat dissipation and reduces temperature rise. The design of the rotor slot number must be matched with the stator slot number. Improper slot matching may lead to motor failure to start, excessive vibration and noise, and additional losses and torque. In this embodiment, the motor uses 54 slots and 42 poles. That is, the number of stator slots 111 on the stator core 11 is 54. The number of magnet groups 22 on the outer wall of the rotor core 21 is 42, forming 42 poles. Please refer to... Figure 4-5 , Figure 4 The waveform height of the 54-slot, 42-pole magnetomotive force is stacked. Figure 5 The histogram of the coefficients for a 54-slot, 42-pole winding is shown below. Figure 4 As can be seen, the combination of 54 slots and 42 poles effectively avoids end-circulating heat generation; from Figure 5 As can be seen, the 54-slot, 42-pole configuration results in a high winding coefficient.

[0064] In this embodiment, one pole consists of 5 magnets. In other embodiments, one pole may consist of 3, 4, or 6 to 10 magnets. By segmenting the magnets, the eddy current losses and heat generation of the magnets during motor operation can be effectively reduced. In this embodiment, N52UH neodymium iron boron magnets are used. The coercivity and squareness coefficient Hk / Hcj information of the N52UH neodymium iron boron magnets are as follows... Figure 7 As shown. From Figure 7 As can be seen from this, N52UH neodymium iron boron magnets can be used for a long time within the range of 180℃, with a square coefficient greater than 0.95 and a remanence Br of 1.46T.

[0065] To ensure sufficient starting torque, the current density cannot be too low, while excessive current density will increase slip, rotor resistance loss, efficiency, and heat generation. In this embodiment, winding 12 is constructed using a concentrated winding method and is a double winding. It is wound with 240℃ temperature-resistant, corona-resistant polyimide copper flat wire, with two wires wound in parallel for four turns. By combining the concentrated winding with the double winding method, the end height can be effectively reduced, thereby reducing copper loss and heat generation during motor operation. In addition, the double winding method can also increase the motor's safety margin, allowing the other winding to continue operating even when one winding is disconnected.

[0066] The stator slot 111 has a fill rate of 89% (including insulating paper). If the slot fill rate is too high, the winding will not be able to fit into the slot; if the slot fill rate is too low, the slot utilization rate will be too low, which is not conducive to the heat dissipation of the winding. In this embodiment, the fill rate of the stator slot 111 is designed to be 89%, which, together with the subsequent design of the stator slot 111 dimensions, ensures both the feasibility of winding copper flat wire and good heat dissipation performance.

[0067] The insulating paper 113 inside the stator slot 111 is made of 0.25mm thick DuPont Nomex T410 aramid insulating paper. The gaps of the winding 12 in the stator slot 111 are filled with insulating varnish 114, which is made of VX4201 unsaturated polyesterimide resin as the slot filling material.

[0068] Rotor 2 is skewed, specifically using a four-segment skew design with angles of -0.5626°, -0.1875°, 0.1875°, and 0.5625°. By employing segmented skewed design with set angles on the rotor, the motor's torque ripple and cogging torque are reduced; motor vibration and noise are also reduced.

[0069] In this embodiment, the stator core 11 and rotor core 21 are made of Hperco50 iron-cobalt-vanadium alloy sheets with a thickness of 0.15mm, laminated together. The stack height of the stator core 11 and rotor core 21 is 65mm, and the stacking factor is 0.97. Please refer to... Figure 6 , Figure 6 The BH curve of Hperco50 iron-cobalt-vanadium alloy material shows the relationship between magnetic induction and magnetic field strength during the magnetization process. As can be seen from the figure, the inflection point is greater than 2.3T and the saturation point is greater than 2.5T.

[0070] The heat sink 3 includes a heat sink ring 31 and a plurality of heat sink fins 32 arranged in a ring at equal angles along the outer periphery of the heat sink ring 31. The heat sink ring 31 is tightly wrapped around the outer periphery of the stator core 11.

[0071] To verify that the motor designed in this embodiment has excellent performance, thermal simulation tests and electromagnetic simulation tests were conducted on the motor, as follows:

[0072] (1) Simulation condition input

[0073] Electromagnetic simulation input conditions: DC bus voltage range 600~710V, DC bus current 0~190A(rms), maximum motor speed 2200rpm; the electromagnetic simulation software is Ansys Motor-CAD, which uses Maxwell's tensor method to perform electromagnetic simulation of the motor. The number of air gap layers is 4, the scanning angle is 360°, and the segment lengths of electromagnetic wires, stator, and rotor are the system defaults. The electromagnetic performance of the motor was simulated under a DC bus voltage of 600V and a DC bus current of 150A (rated power) and 190A (peak power).

[0074] Thermal simulation input conditions: maximum ambient temperature 55℃, air volume exchanging heat with the motor surface per minute is 2000 liters; at rated power: electrical load, electrical density, and thermal load: 67.44 A / mm², 12.78 A / mm², respectively. 2 861.89A 2 / mm 3 After thermal equilibrium, the highest winding temperature is 155.1℃, and the magnet temperature is 106.2℃; under peak power: electrical load, electrical density, and thermal equivalent are 85.43 A / mm and 16.19 A / mm, respectively. 2 1383.07A 2 / mm 3 After 120 seconds, the highest winding temperature was 157.7℃, and after 180 seconds it was 177.6℃.

[0075] (2) Simulation process

[0076] (21) Electromagnetic simulation at peak power:

[0077] The current input is 190A, the voltage input is 600V, the lead angle is set to 10°, the winding copper wire temperature is set to 173℃, the magnet temperature is set to 157℃, the shaft temperature is set to 105℃, the maximum speed is 2200rpm, the iron loss coefficient Ka is 1.7, and the electrical load, electrical density, and thermal equivalent are 85.43A / mm, 16.19A / mm, and 16.19A / mm, respectively. 2 1383.07A 2 / mm 3 .

[0078] Electromagnetic simulation data at peak power, such as Figure 8 As shown in the figure. The simulation results show that the maximum torque is 553.56 Nm at 1923 rpm, the highest efficiency is 96.32%, and the total loss is 4868.7 W.

[0079] Magnetic induction density cloud map at peak power as shown Figure 9 As shown; the magnetic induction intensity of each part is as follows Figure 10 As shown. Simulation results show that the maximum magnetic flux density is located in the tooth section, with a maximum of 2.425T, the average magnetic flux density in the air gap is 0.8012T, and the maximum magnetic flux density in the air gap is 1.287T.

[0080] The loss distribution obtained from the simulation at peak power is as follows: Figure 11 As shown. Simulation results show: peak shaft power 109.80KW, 553.56Nm@1923rpm, efficiency 96.32% (winding temperature 173℃, magnet temperature 157℃).

[0081] The TS curve obtained from the simulation at peak power is as follows: Figure 12 As shown in the figure. The simulation results show that the maximum torque of 553.56 Nm is met within the speed range of 1923 rpm, and the torque is still greater than 540 Nm at 2200 rpm.

[0082] The efficiency MAP obtained from the simulation at peak power is shown below. Figure 13 As shown in the figure, the simulation results (MAP) for peak power show that the motor efficiency is greater than 96% in the speed range of 1000rpm to 2200rpm.

[0083] (22) Thermal simulation at peak power:

[0084] When the aircraft climbs vertically at a rate of 3 meters per second, the motor casing displaces 2260.8 liters of air per minute. Taking 0.88 times the altitude at sea level at 3000 meters as an example, and considering all factors, the amount of air in contact with the motor casing per minute is estimated to be 2000 liters.

[0085] Thermal simulation results at peak power are as follows Figure 14 As shown in the figure. The simulation results show that: under an external environment of 55℃ and an altitude of 3000 meters, the winding temperature is about 157.7℃ after 120 seconds and 177.6℃ after 180 seconds. The motor reaches thermal equilibrium within 10 minutes of operation, with the winding temperature at about 211.5℃ and the magnet temperature at about 126.3℃. There is no winding burnout or magnet demagnetization, indicating a steady state.

[0086] (21) Electromagnetic simulation at rated power:

[0087] The current input is 150A, the voltage input is 600V, the lead angle is set to 10°, the winding copper wire temperature is set to 135℃, the magnet temperature is set to 130℃, the shaft temperature is set to 85℃, the maximum speed is 2200rpm, the iron loss coefficient Ka is 1.7, and the electrical load, electrical density, and thermal load are 67.44A / mm, 12.78A / mm, and 12.78A / mm, respectively. 2 861.89A 2 / mm 3 .

[0088] Electromagnetic simulation data at rated power, such as Figure 15 As shown in the figure. The simulation results show that the maximum torque is 455.09 Nm @ 2018.8 rpm, the highest efficiency is 97.05% (winding temperature 135℃, magnet temperature 130℃), and the total loss is 3190.1 W.

[0089] Magnetic induction density cloud diagram at rated power as shown Figure 16 As shown, the magnetic induction intensity of each part is as follows Figure 17 As shown. Simulation results show that the maximum magnetic flux density is located in the tooth section, with a maximum of 2.365T, the average magnetic flux density in the air gap is 0.8252T, and the maximum magnetic flux density in the air gap is 1.297T.

[0090] The loss distribution obtained from the simulation at rated power is as follows: Figure 18 As shown.

[0091] The TS curve obtained by simulation under rated power is as follows: Figure 19 As shown in the figure. The simulation results show that the maximum torque of 455.09 Nm is met within the speed range of 2018.8 rpm, and the torque is still greater than 450 Nm at 2200 rpm.

[0092] The efficiency MAP obtained from the simulation at rated power is shown below. Figure 20 As shown in the figure, the simulation results (MAP) for peak power efficiency show that the motor efficiency is greater than 96.5% in the speed range of 1000rpm to 2200rpm.

[0093] (23) Thermal simulation at rated power:

[0094] When the aircraft climbs vertically at a rate of 3 meters per second, the motor casing displaces 2260.8 liters of air per minute. Taking 0.88 times the altitude at sea level at 3000 meters as an example, and considering all factors, the amount of air in contact with the motor casing per minute is estimated to be 2000 liters.

[0095] Thermal simulation results at rated power are as follows Figure 21 As shown. Simulation results show that in an external environment of 55℃ and at an altitude of 3000 meters, the winding temperature is approximately 118.3℃ after 120 seconds and 144.8℃ after 300 seconds. The motor reaches thermal equilibrium within 15 minutes of operation, with the winding temperature at 155.1℃ and the magnet temperature at approximately 106.2℃. There is no winding burnout or magnet demagnetization.

[0096] (24) Maximum torque simulation:

[0097] Maximum torque electromagnetic simulation input conditions: current input 190A, voltage input 600V, lead angle set to 10°, winding copper wire temperature set to 20℃, magnet set to 20℃, shaft temperature set to 20℃, maximum speed 2200rpm.

[0098] The simulation results of the maximum torque are as follows: Figure 22 As shown in the figure. The simulation results show that the maximum torque is 650.89 Nm, the highest efficiency is 97.48%, and the total loss is 3879.4 W.

[0099] The magnetic flux density cloud diagram at maximum torque is shown below. Figure 23 As shown, the magnetic induction density of each part is as follows: Figure 24 As shown in the figure. Simulation results show that at maximum torque output, the highest magnetic induction intensity is 2.505, the maximum air gap magnetic flux density is 1.425T, and the average is 0.9505T.

[0100] (25) Rotor strength simulation. The simulation results are as follows: Figure 25-26 As shown, the results indicate that at a rotational speed of 2200 rpm, the rotor did not reach the yield point of 393 MPa for Hperco50 material, and the actual maximum stress was less than 18 MPa.

[0101] (26) NVH performance testing. Simulation results of cogging torque and torque pulsation are as follows: Figure 27 As shown in the figure. Simulation results show that with a 4-segment rotor skewed pole treatment, the torque ripple is less than 2%, and the cogging torque is only 1.305 Nm.

[0102] (27) Time-dominant stator radial harmonic 2D spectrum, as shown Figure 28 As shown in the figure, the results indicate that the wave peaks are rounded and smooth, effectively avoiding the whine problem during motor operation.

[0103] (28) Campbell's sound pressure field spectrum, as shown Figure 29 As shown in the figure. The results show that: spatial orders 6 and 12 correspond to NVH abnormalities in the frequency range of 1500-2300 Hz, and corresponding to speeds above 4500 rpm (2200 rpm, 733.33 Hz).

[0104] (29) Stator radial stress data, such as Figure 30 As shown.

[0105] (30) The phase current curve, peak terminal voltage, and back electromotive force (EMF) curve are respectively as follows: Figures 31-33 As shown in the figure, the peak voltage is 268.66A, the peak terminal voltage is 684.1V, and the peak back EMF is 589.8V at 2200rpm.

[0106] Based on the above simulation experiments, the motor parameter list for this embodiment is shown in the table below.

[0107]

[0108]

[0109] For those skilled in the art, various modifications and improvements can be made without departing from the inventive concept of this invention, and these all fall within the protection scope of this invention.

Claims

1. A high torque density electric motor for an aircraft, comprising a stator and a rotor disposed inside the stator; the stator comprising a stator core and windings, wherein the inner wall of the stator core is provided with a plurality of stator slots at equal angles along its circumference, and each stator slot is provided with the windings; the rotor comprising a rotor core and a set of magnets, wherein the outer wall of the rotor core is provided with a plurality of receiving cavities at equal angles along its circumference, and each receiving cavity is provided with a set of magnets; characterized in that: The stator has 54 slots and the magnet assembly has 42, forming a 54-slot, 42-pole configuration. The winding adopts a concentrated double winding form; The stator core has an outer diameter D1 = 355 mm and an inner diameter D2 = 308.6 mm. The outer diameter of the rotor core is D3=306mm and the inner diameter is D4=286mm. The air gap L1 between the stator and the rotor is 1.3 mm; The stator slot has a slot opening width L2=3.5mm, a slot depth L3=18.7mm, a slot tip angle θ1=28°, and the stator slot is a parallel slot with a parallel slot width L4=9.4mm. The rotor adopts a four-segment skewed pole treatment, with skewed pole treatment angles of -0.5626°, -0.1875°, 0.1875°, and 0.5625° respectively. The stator core and the rotor core are made of Hperco50 iron-cobalt-vanadium alloy sheets with a thickness of 0.15mm, and the stack height of the stator core and the rotor core is 65mm, with a stacking factor of 0.

97.

2. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: Stator teeth are formed between two adjacent stator slots, and the air gap is formed between the inner wall of the stator teeth and the outer wall of the magnet assembly.

3. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: The main cross-sectional shape of the stator slot is rectangular, and an inwardly recessed slot is formed on the side near the inner wall of the stator core, with the inwardly recessed profile being a straight line; the slot tip angle is the inclination angle of the inwardly recessed profile.

4. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: The magnets in the magnet assembly have a thickness L5 = 5.2 mm and are inserted into the receiving cavity in a radial insertion manner.

5. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: The magnet is an N52UH neodymium iron boron magnet with a remanence of 1.46T and a maximum operating temperature of 180℃.

6. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: The group of magnets forms a pole, and the pole includes 6 to 10 magnets.

7. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: The winding is made of 240℃ temperature-resistant, corona-resistant polyimide copper flat wire.

8. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: The stator slot fill factor is 89.4%.

9. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: The stator slot is provided with insulating paper that isolates the inner wall of the stator slot from the winding; the insulating paper is 0.25mm thick DuPont Nomex T410 aramid insulating paper; the gap of the winding in the stator slot is filled with insulating varnish, which is VX4201 unsaturated polyesterimide resin.

10. The high torque density electric motor for electric aircraft as described in claim 1, characterized in that: The stator is provided with a heat dissipation component on its outer periphery; the heat dissipation component includes a heat dissipation ring and a plurality of heat dissipation fins arranged in a ring at equal angles along the outer periphery of the heat dissipation ring; the heat dissipation ring is tightly wrapped around the outer periphery of the stator core.

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